Three-Dimensional Printing of Ultrasoft Silicone with a Functional Stiffness Gradient

A methodology for three-dimensionally printing ultrasoft silicone with a functional stiffness gradient is presented. Ultraviolet-cure silicone was deposited via two independently controlled extruders into a thixotropic, gel-like, silicone oil-based support matrix. Each extruder contained a different liquid silicone formulation. The extrusion rates were independently varied during printing such that the combined selectively deposited material contained different ratios of the two silicones, resulting in localized control of material stiffness. Tests to validate the process are reported, including tensile testing of homogeneous cubic specimens to quantify the range of material stiffness that could be printed, indentation testing of cuboid specimens to characterize printed stiffness gradients, and vibratory testing of synthetic multilayer vocal fold (VF) models to demonstrate that the method may be applied to the fabrication of biomechanical models for voice production research. The cubic specimens exhibited linear stress–strain data with tensile elasticity modulus values between 1.11 and 27.1 kPa, more than a factor of 20 in stiffness variation. The cuboid specimens exhibited material variations that were visually recognizable and quantifiable via indentation testing. The VF models withstood rigorous phonatory flow-induced vibration and exhibited vibratory characteristics comparable to those of previous models. Overall, while process refinements are needed, the results of these tests demonstrate the ability to print ultrasoft silicone with stiffness gradients.

another.This is often done by either retracting the current material from the extruder and then priming the new material in the same extruder using a purge tower, or by switching between multiple extruders with different materials.Material changes in standard multi-color FFF 3D printing are time consuming, and as a result, slicers are typically designed to minimize the number of material changes.While this may save time, it requires a significant amount of travel movement where no material is being deposited.Because FFF 3D printing with thermoplastics is not susceptible to unwanted flow, the additional travel movement required to minimize material changes is inconsequential.However, for silicone 3D printing, the unwanted flow caused by excess travel due to material changes can cause unwanted deformities in the final print.Excess needle movement caused by infill-patterns is the second example of unwanted flow.
Standard 3D-printing slicing software typically offers a variety of print infill-patterns that can be selected.For single-material 3D-printed synthetic vocal fold models, the simple rectilinear pattern with 45° infill-orientation, as shown in Fig. S.1, was found to yield the best printing results; 24 however, these same results did not translate well to multi-material 3D printing. 28In the latter case, the infill-pattern is applied to each material section as shown in The custom slicing software was designed to overcome the challenges of unwanted flow.
The software optimized the print order to reduce travel movement (Fig The slicing software then "sliced" the print layer-by-layer creating x-, y-, and z-translation g-code commands for the 3D printer.The layer height was equal to the needle inner diameter (0.21 mm), and the path width of the continuous rectilinear print pattern was equal to a multiple of the needle inner diameter and the percent infill (60%).Other settings that were implemented during the slicing phase of the software include print speed, extruder retraction distance, infill angle, syringe diameter, needle outer diameter, and needle offset.
After the print was "sliced," the part in which each path was located was determined, and the corresponding ratio of extruder A and extruder B was assigned.This ratio was used to calculate the A-and B-extruder g-code commands.Additionally, each path was checked to see if a material change was required, and necessary g-code was added accordingly.Finally, the g-code for each layer was compiled and exported for the entire print.
The software provided a layer-by-layer animation of the print as shown in Fig

Finite Element Model
It was expected that compression testing near the cuboid edges as well as near interfaces of material sections with different stiffnesses would affect compression results.Therefore, for comparison and to confirm the effects of edges and material interfaces on the stiffness profiles, finite element analysis (FEA) of each cuboid compression test was performed.An FEA model of the 10×10×33 mm cuboid was created using ANSYS ADPL (Canonsburg, PA, see Young 29 for code).Each material section of the cuboid was modeled using 1 mm high-order SOLID187 3D elements with a bonded boundary condition between each material section.The section materials were modeled using a linear elastic material with a Poisson's ratio of 0.49, and the bottom of the cuboid was fixed in all directions.After the initial meshing, the elements in a hemisphere around the region of indentation were refined to improve the FEA results, with the refinement increasing closer to the contact area as shown in Fig. S.6.The first refinement of the elements was within a 9.6 mm diameter sphere of the compression testing location.The second refinement was within a smaller hemisphere with a diameter 6.4 mm.The last refinement was on the surface of the cuboid, using the same refinement settings, to provide a high density of nodes in a circular pattern with a 3.4 mm diameter.
The initial FEA analyses modeled the indenter as a cylinder and contact pairing with the cuboid.An alternative approach was tested in which the indenter was modeled as a 3 mm diameter circle inscribed on the top of the cuboid with the center being located at the desired testing location.
In both approaches, the cylinder or circle was displaced 2 mm into the cuboid over 40 load steps (i.e., 0.05 mm penetration at each load step) and fixed in all other remaining dimensions.The differences in initial cuboid stiffness results were found to be negligible; however, the computational cost of the cylinder method was higher.Therefore, the final FEA analyses were completed using an inscribed circle rather than a cylinder.The initial FEA analyses were solved with nlgeom activated in ANSYS to account for large deflections; however, above load step 17 (0.85 mm) the solution frequently became nonconvergent.Of the load steps that did solve, the cuboid stiffnesses with nlgeom activated and deactivated for the n = 2 and n = 11 cuboids were a maximum of 3.62% and 3.83% different, respectively, with averages of 1.44% and 1.35%, respectively.With the relatively small average differences between nlgeom activated and deactivated, as well as the linearity of the experimental force-displacement data shown in Young, 29 the material for the FEA model was assumed to be linear-elastic and the final FEA analyses were completed with nlgeom deactivated.
For the initial FEA analyses, each sections' material model was assigned a uniform modulus value that corresponded to the target modulus values of the printed cuboids (i.e., 1.5 kPa and 12 kPa for the two sections of the n = 2 cuboid, and 1.50 kPa, 2.55 kPa, 3.60 kPa, 4.65 kPa, 5.70 kPa, 6.75 kPa, 7.80 kPa, 8.85 kPa, 9.90 kPa, 10.95 kPa, and 12 kPa for the 11 sections of the n = 11 cuboid).The FEA stiffness results were then analyzed and compared to the experimental results.The FEA modulus values were then tuned to adjust the stiffness results to more closely match the experimental compression data and the FEA analyses were repeated.The modulus values were tuned by scaling the initial FEA modulus input values by the ratio of the desired FEA stiffness over the actual FEA stiffness at points near the ends of the cuboids.For the n = 11 cuboid, the 4 th and 22 nd data points were selected, and for the n = 2 cuboids, the 4 th and 18 th data points were selected.The primary purpose behind performing FEA analyses was to validate the trends (i.e., stiffness gradients, edge effects, and material interfaces) that were present in the experimental data, therefore, stiffness tuning was deemed to be an acceptable approach to better compare the trends.Final FEA modulus values are listed in Table S.1, and a comparison between pre-and posttuned stiffness plots are shown in Fig. 6.
FIG. S.1.Print order for sections of a multi-material layer.Each color represents a different material.(a) All of one material is printed before switching to the next material.(b) The layer is printed from left to right continuously to minimize travel movement.
FIG. S.2.Two infill-patterns.(a) Rectilinear pattern with 45° infill orientation applied to each material section causing the needles to turn around and create separation at each material section.(b) Rectilinear infill-pattern applied across the entire print layer creating a continuous infill-pattern with material switching at the interface of different sections.
FIG. S.4.Illustration of model fabrication process.
FIG. S.5.Image of the slicing software graphical user interface showing the g-code output (left panel), layer-by-layer animation (center panel), and extruder values (right panel).The interface allows the user to easily navigate between layers and g-code lines to verify correct g-code output.
FIG. S.6.(Left) Perspective view of the meshed n = 2 cuboid.Refined circular mesh area is located where the indenter is plunged into the cuboid in this instance.(Right) Perspective view of the deformed cuboid with 2 mm indentation.The same meshing and indentation procedure was followed the for the n = 11 cuboid.

Table S .
1. Post-tuned lower and upper modulus values for the cuboid FEA models.The initial (pre-tuned) lower and upper modulus values were 1.5 and 12 kPa, respectively.